Method And Apparatus For Calculating Insertion Indeces For A Modular Multilevel Converter

ABSTRACT

A method for calculating insertion indices for a phase leg of a DC to AC modular multilevel converter. Each phase leg includes two serially connected arms, wherein each arm includes a number of submodules, wherein each submodule can be in a bypass state or a voltage insert mode. The insertion index includes data representing the portion of available submodules that should be in the voltage insert mode. The method includes the steps of: calculating a desired arm voltage for an upper arm connected to the upper DC source common bar and a lower arm connected to the lower DC source common bar, obtaining values representing actual total arm voltages in the upper arm and lower arm, respectively, and calculating modulation indices for the upper and lower arm, respectively, using the respective desired arm voltage and the respective value representing the total actual arm voltage. A corresponding apparatus is also presented.

CROSS-REFERENCE TO RELATED APPLICATIONS

The present application is a continuation of pending International patent application PCT/EP2010/062923 filed on Sep. 3, 2010 which designates the United States and claims the benefit under 35 U.S.C. §119 (e) of the U.S. Provisional Patent Application Ser. No. 61/239,859 filed on Sep. 4, 2009. The content of all prior applications is incorporated herein by reference.

FIELD OF THE INVENTION

The invention relates to the calculation of insertion indices comprising data representing the portion of available submodules of a modular multilevel converter that should be in voltage insert mode.

BACKGROUND OF THE INVENTION

The concept Modular Multilevel Converter (M2C) denotes a class of Voltage Source Converter (VSC). It has one or several phase legs connected in parallel between two DC bars, a positive DC+ and a negative DC−. Each phase leg consists of two series-connected converter arms. The connection point between the converter arms constitutes an AC terminal for the leg.

Each arm consists of a number (N) of submodules. Each submodule has two terminals. Using these terminals the submodules in each arm are series-connected so that they form a string. The end terminals of the string constitute the connection terminals of the arm. By controlling individual modules in each arm, a voltage corresponding to the accumulation of insertion voltages can be provided on the AC terminal.

Such a converter is known from DE10103031. In this document, a method to equalize the voltages in the submodule capacitors within the arm is briefly described. For each arm, a modulator determines when the number of inserted submodules shall change. The principle for equalizing is that, at each instant when a change of the number of inserted submodules is commanded, a selection mechanism chooses the submodule to be inserted or bypassed depending on the actual current direction in the arm (charging or discharging) and the corresponding available submodules in the arm (bypassed highest voltage/bypassed lowest voltage/inserted highest voltage/inserted lowest voltage). Such a selection mechanism aims to achieve that the DC voltage across the DC capacitors in the submodules are equal, u_(C,SM)(t).

A problem with the prior art is the presence of a circulation current going through the legs between the DC terminals.

SUMMARY OF THE INVENTION

An object of the invention is to reduce circulation currents going through the legs between the DC terminals.

A first aspect is a method for calculating insertion indices for a phase leg of a DC to AC modular multilevel converter, the converter comprising one phase leg between upper and lower DC source common bars for each phase, each phase leg comprising two serially connected arms, wherein an AC output for each phase leg is connected between its two serially connected arms. Each arm comprises a number of submodules, wherein each submodule can be in a bypass state or a voltage insert mode, the insertion index comprising data representing the portion of available submodules that should be in the voltage insert mode for a particular arm. The method comprises the steps of: calculating a desired arm voltage for an upper arm connected to the upper DC source common bar and a lower arm connected to the lower DC source common bar, obtaining values representing actual total arm voltages in the upper arm and lower arm, respectively, and calculating insertion indices for the upper and lower arm, respectively, using the respective desired arm voltage and the respective value representing the total actual arm voltage.

The step of calculating desired arm voltages for a phase leg may comprise calculating

u _(CU)(t)=u _(D)/2−e _(V)(t)−u _(diff)(t)

for the upper arm, and calculating

u _(CL)(t)=u _(D)/2+e _(V)(t)−u _(diff)(t)

for the lower arm, where u_(CL)(t) represents upper arm voltage, u_(D) represents a voltage between the upper and lower DC source common bars, e_(V)(t) represents a reference inner AC output voltage and u_(diff)(t) represents a control voltage to control a current passing through the whole phase leg.

The step of calculating a desired arm voltage may comprise calculating

u _(diff)(t)=u _(diff1)(t)+u _(diff2)(t)

where u_(diff1)(t) represents a voltage obtained by summing energy in the arms of the leg and u_(diff2)(t) represents a voltage obtained by calculating a difference in energy between the arms of the leg.

The step of obtaining a value representing actual arm voltage may comprise calculating

u _(diff2)(t)=û _(diff2) cos(ω₁ t+ψ)

where û_(diff2) represents an error between total upper arm energy and total lower arm energy, ω₁ represents the angular velocity of the network frequency and iv represents the angle given by ψ=<(R+jω₁L) where R represents the resistance of the converter arm and L represents the inductance of the converter arm.

The step of obtaining values representing actual arm voltages may comprise: calculating u_(CU) ^(Σ)(t), actual voltage for the upper arm, using C_(arm), capacitance for the arm, î_(diff0), DC current passing through the two serially connected arms of the phase leg, W_(CU) ^(Σ)(t), desired average energy in the upper arm, ê_(V), amplitude of reference for the inner AC output voltage, î_(V), amplitude of AC output current, φ, a phase difference between i_(V)(t) and e_(V)(t), DC current circulating through the two series-connected arms, and calculating u_(CL) ^(Σ)(t), actual voltage for the lower arm, using C_(arm), capacitance for the arm, î_(diff0), DC current passing through the two serially connected arms of the phase leg, W_(CL) ^(Σ)(t), desired average energy in the lower arm, ê_(V), amplitude of reference for inner AC output voltage, î_(V), amplitude of AC output current φ, a phase difference between i_(V)(t) and e_(V)(t), î_(diff0), DC current circulating through the two series-connected arms.

The step of obtaining a value representing actual arm voltage may comprise calculating

${\hat{i}}_{diff0} = \frac{{\hat{e}}_{v}{\hat{i}}_{v}\cos \; \phi}{u_{D} + \sqrt{u_{D}^{2} - {4R{\hat{e}}_{v}{\hat{i}}_{v}\cos \; \phi}}}$

where φ represents a phase difference between i_(V)(t) and e_(V)(t), u_(D) represents a voltage between the upper and lower DC source common bars and R represents the resistance of the converter arm.

The step of obtaining a value representing actual arm voltage may comprise calculating

${u_{CU}^{\Sigma}(t)} = \sqrt{\frac{2{W_{CU}^{\Sigma}(t)}}{C_{arm}}}$

where W_(CU) ^(Σ)(t) represents instantaneous energy in the upper arm and is calculated as follows:

${W_{CU}^{\Sigma}(t)} = {W_{{CU}\; 0}^{\Sigma} - {\frac{{\hat{e}}_{v}{\hat{i}}_{{diff}\; 0}}{\omega_{1}}\sin \; \omega_{1}t} + {\left( {\frac{u_{D}}{2} - {R\; {\hat{i}}_{{diff}\; 0}}} \right)\frac{{\hat{i}}_{V}}{2\omega_{1}}{\sin \left( {{\omega_{1}t} + \phi} \right)}} - {\frac{{\hat{e}}_{V}{\hat{i}}_{V}}{8\omega_{1}}{\sin \left( {{2\omega_{1}t} + \phi} \right)}}}$

and calculating

${u_{CL}^{\Sigma}(t)} = \sqrt{\frac{2{W_{CL}^{\Sigma}(t)}}{C_{arm}}}$

where W_(CL) ^(Σ)(t) represents instantaneous energy in the lower arm and is calculated as follows:

${W_{CL}^{\Sigma}(t)} = {W_{{CL}\; 0}^{\Sigma} - {\frac{{\hat{e}}_{v}{\hat{i}}_{{diff}\; 0}}{\omega_{1}}\sin \; \omega_{1}t} - {\left( {\frac{u_{D}}{2} - {R\; {\hat{i}}_{{diff}\; 0}}} \right)\frac{{\hat{i}}_{V}}{2\omega_{1}}{\sin \left( {{\omega_{1}t} + \phi} \right)}} - {\frac{{\hat{e}}_{V}{\hat{i}}_{V}}{8\omega_{1}}{\sin \left( {{2\omega_{1}t} + \phi} \right)}}}$

where ω₁ represents the angular velocity of the network frequency, u_(D) represents a voltage between the upper and lower DC source common bars, R represents the resistance of the converter arm.

The step of obtaining a value representing actual arm voltage may comprise measuring voltages of the submodules of the arm and summing these measured voltages.

The insertion index may comprise data representing a direction of the inserted voltage.

A second aspect is an apparatus for calculating insertion indices for a phase leg of a DC to AC modular multilevel converter, the converter comprising one phase leg between upper and lower DC source common bars for each phase, each phase leg comprising two serially connected arms, wherein an AC output for each phase leg is connected between its two serially connected arms, wherein each arm comprises a number of submodules. Each submodule can be in a bypass state or a voltage insert mode, the insertion index comprising data representing the portion of available submodules that should be in the voltage insert mode for a particular arm. The apparatus comprises a controller arranged to calculate a desired arm voltage for an upper arm connected to the upper DC source common bar and a lower arm connected to the lower DC source common bar, to obtain values representing actual total arm voltages in the upper arm and lower arm, respectively, and to calculate modulation indices for the upper and lower arm, respectively, using the respective desired arm voltage and the respective value representing the total actual arm voltage.

Generally, all terms used in the application are to be interpreted according to their ordinary meaning in the technical field, unless explicitly defined otherwise herein. All references to “a/an/the element, apparatus, component, means, step, etc.” are to be interpreted openly as referring to at least one instance of the element, apparatus, component, means, step, etc., unless explicitly stated otherwise. The steps of any method disclosed herein do not have to be performed in the exact order disclosed, unless explicitly stated.

BRIEF DESCRIPTION OF THE DRAWINGS

The invention is now described, by way of example, with reference to the accompanying drawings, in which:

FIG. 1 is a schematic diagram of phase legs and arms;

FIG. 2 is a schematic circuit model of a phase leg of FIG. 1;

FIG. 3 is a Nichols plot for an open loop transfer function;

FIG. 4 is a Nichols plot for an open loop transfer function when a PID controller is used;

FIG. 5 is a Nichols plot for an open loop transfer function when a PID controller, time delay and notch filter is used;

FIG. 6 is a graph showing a simulation result at a step in the reference for the total energy in the converter leg;

FIG. 7 is a graph showing the simulation result when the current changes from 0.1 pu to 1.0 pu in the converter leg; and

FIG. 8 shows a Nichols plot for a balance controller according to another embodiment.

DETAILED DESCRIPTION OF THE INVENTION

The invention will now be described more fully hereinafter with reference to the accompanying drawings, in which certain embodiments of the invention are shown. This invention may, however, be embodied in many different forms and should not be construed as limited to the embodiments set forth herein; rather, these embodiments are provided by way of example so that this disclosure will be thorough and complete, and will fully convey the scope of the invention to those skilled in the art. Like numbers refer to like elements throughout the description.

In the description in the following, continuous variables are used, corresponding to the simplifying assumption that the arms have infinite number of submodules that are switched with infinite switching frequency.

FIG. 1 shows an M2C (Modular Multilevel Converter) having a phase leg 7 that comprising an upper arm 5 and a lower arm 6. Each arm 5, 6 comprising a number of serially connected submodules 9. Each submodule 9 comprises a switchable capacitor. An AC output 8 is connected between the upper and lower arms 5, 6. Although only one phase leg 7 is shown here, the M2C comprises one phase leg 7 for each phase, i.e. three phase legs 7 for a three phase system, where each phase leg 7 comprises upper and lower arms 5, 6 comprising submodules 9.

An upper DC source common bar (in this case DC+) and a lower DC source common bar (in this case DC−) for each phase is provided. It is to be noted that the upper and lower DC source common bars can switch polarity.

Ideally the capacitors keep a constant DC voltage and the AC terminal voltage is controlled by varying the number of inserted modules in the upper and lower arms. If the voltage between the DC bars is constant this obviously requires that, in average, the total number of inserted modules in the two arms remain constant. The arm inductors however will limit the rate of change of the arm currents, making it possible to accept minor short deviations from this condition.

Now once the context is presented, define the insertion index, n_(x)(t), for the arm x to be the ratio between the inserted number of submodules and the total number of available submodules in the arm. The arm voltage then becomes

u _(Cx)(t)=n _(x)(t)u _(C,SM)(t)  (1)

In a simple approach the number of inserted modules in each arm can be generated by the modulator much in the same way as in PWM modulation for conventional VSCs. Then, in order to generate an inner AC voltage with amplitude ê_(V) the insertion indices for the upper and the lower arms become

$\begin{matrix} {{{u_{CU}(t)} = {{{n_{U}(t)}u_{D}\mspace{14mu} {u_{CL}(t)}} = {{n_{L}(t)}u_{D}}}}{{n_{U}(t)} = {{\frac{1 - {\hat{m}\; \cos \; \omega \; t}}{2}\mspace{20mu} {n_{L}(t)}} = \frac{1 - {\hat{m}\; \cos \; \omega \; t}}{2}}}{\hat{m} = \frac{{\hat{e}}_{V}}{\frac{u_{D}}{2}}}} & (2) \end{matrix}$

When the simple modulation approached described above is used and the converter is loaded on its AC side the desired waveform will be distorted due to the ripple in the capacitor voltages that will be created when the load current passes through the converter arms. Specifically a strong second harmonic current will circulate through the converter leg and the DC side and/or the neighbor phases. This undesired second harmonic current increases the peak of the arms currents and causes extra losses in the converter arms.

The problem can be solved by generating the insertion indices for the arms, n_(U) and n_(L), in other ways. Such methods would aim to:

-   -   eliminate and/or control the harmonic current in the converter         arms     -   for each arm control the total energy stored in all capacitors         in that arm which is equivalent to control the total voltage of         all capacitors in the arm     -   thereby control the total energy stored in the phase leg as well         as the balance between the upper and the lower arms in the phase         leg

According to the invention the insertion indices n_(U)(t) and n_(L)(t) for the converter arms are being derived in real-time according to the following procedure

-   -   the reference for the converter inner voltage relative the         midpoint of the DC link is given in the form e_(V)(t)=ê_(V) cos         ω₁t; this reference typically is delivered by an AC side         controller operating on AC quantities like output voltage,         current or flux; the converter circuit parameters like arm         resistance and inductance may be used by the controller     -   the desired arm voltages u_(CU)(t) and u_(CL)(t) are calculated         as

$\begin{matrix} {{u_{CU}(t)} = {{\frac{u_{D}}{2} - {e_{V}(t)} - {{u_{diff}(t)}\mspace{14mu} {u_{CL}(t)}}} = {\frac{u_{D}}{2} + {e_{V}(t)} - {u_{diff}(t)}}}} & (3) \end{matrix}$

-   -   where u_(D) is the voltage between the DC rails and u_(diff)(t)         is a control voltage that is created by the control system that         will be described later in the this memo     -   the total capacitor voltages, u_(CU) ^(Σ)(t) and u_(CL) ^(Σ)(t),         of all capacitors in the upper and lower arms respectively, are         measured or estimated as will be described later in this memo     -   the insertion indices are calculated as

$\begin{matrix} {{n_{U}(t)} = {{\frac{u_{CU}(t)}{u_{CU}^{\Sigma}(t)}\mspace{14mu} {n_{L}(t)}} = \frac{u_{CL}(t)}{u_{CL}^{\Sigma}(t)}}} & (4) \end{matrix}$

According to the invention there are two different ways to create the variables u_(diff)(t), u_(CU) ^(Σ)(t) and u_(CL) ^(Σ)(t).

In this approach the sum of the capacitor voltages in each arm, u_(CU) ^(Σ)(t) and u_(CL) ^(Σ)(t), are measured using sensors in the submodules. If the voltage sharing between the modules is assumed to be even the total energies in each arm can be calculated as

$\begin{matrix} {{W_{CU}^{\Sigma}(t)} = {{\frac{C_{arm}}{2}\mspace{14mu} {u_{CU}(t)}^{2}\mspace{14mu} {W_{CL}^{\Sigma}(t)}} = {\frac{C_{arm}}{2}\mspace{14mu} {u_{CL}(t)}^{2}}}} & (5) \end{matrix}$

where C_(arm)=C_(submod)/N. Alternatively the energy in each arm capacitor can be calculated individually and the total energy for each arm then can be created by summing the energies in all submodules in each arm. The voltage reference component u_(diff)(t) is created as the sum of the output signals from two independent controllers u_(diff)(t)=u_(diff1)(t)+u_(diff2)(t).

The first controller has a reference for the total energy in both arms of the phase leg. The response signal is the measured total energy W_(CU) ^(Σ)(t)+W_(CL) ^(Σ)(t) which may be filtered using e.g. a notch filter tuned to the frequency 2ω₁ (ω₁ is the network frequency) or any other filter suppressing the same frequency. The error, i.e. the difference between the reference and the response signals, is connected to a controller (normally of type PID) that has the output signal u_(diff1)(t).

The second controller has a reference for the difference between the energies in the arms in the phase leg. This reference typically is zero, meaning that the energy in the arms in the phase leg shall be balanced. The response signal is created as the measured W_(CU) ^(Σ)(t)−W_(CL) ^(Σ)(t), filtered by a notch filter tuned to ω₁ or any other filter suppressing the same frequency. The error is brought to a controller (typically of P type), which has an output signal û_(diff2). The contribution to the total voltage reference u_(diff)(t) is obtained by multiplying û_(diff2) by a sinusoidal time function cos(ω₁t+ψ), which is phase-shifted relative the inner voltage reference by the angle ψ given by ψ=<(R+jω₁L), where R and L are the resistance and inductance respectively in the converter arm. Thus:

u _(diff)(t)=u _(diff1)(t)+û _(diff2) cos(ω₁ t+ψ)  (6)

The first approach to stabilisation of the converter according to the procedure described in this section is described in more detail in Appendix 1.

Remark 1: The reference for the AC side inner voltage may comprise a minor third harmonic voltage component, which is used to increase the available output voltage level in a 3-phase converter. This does not impact significantly on the behaviour described.

Remark 2: A third reference component may be added to the control voltage u_(diff)(t). This component has the purpose of intentionally creating a second harmonic current in the arms in order to increase the available output voltage for loads with certain power factors.

Second Approach, Open-Loop Control

In this approach the AC side current i_(V(t)) is measured. Its fundamental frequency component is extracted with amplitude and phase relative the reference inner voltage e_(V(t)) for the converter. Thus the AC side current can be written as

i _(V)(t)=î _(V) cos(ω₁ t+φ)  (7)

Assuming that the converter shall operate ideally in steady-state, i.e. it shall produce undisturbed AC output voltage and the upper and lower arms shall carry half the AC output current each, it is possible to calculate the ideal derivative of the energies in each arm. The result is

$\begin{matrix} {\frac{W_{CL}^{\Sigma}}{t} = {{{- {\hat{e}}_{V}}{\hat{i}}_{{diff}\; 0}\cos \; \omega_{1}t} + {\left( {\frac{u_{D}}{2} - {R\; {\hat{i}}_{{diff}\; 0}}} \right)\frac{{\hat{i}}_{V}}{2}{\cos \left( {{\omega_{1}t} + \phi} \right)}} - {\frac{{\hat{e}}_{V}{\hat{i}}_{V}}{4}{\cos \left( {{2\omega_{1}t} + \phi} \right)}}}} & (8) \\ {\frac{W_{CL}^{\Sigma}}{t} = {{{+ {\hat{e}}_{V}}{\hat{i}}_{{diff}\; 0}\cos \; \omega_{1}t} - {\left( {\frac{u_{D}}{2} - {R\; {\hat{i}}_{{diff}\; 0}}} \right)\frac{{\hat{i}}_{V}}{2}{\cos \left( {{\omega_{1}t} + \phi} \right)}} - {\frac{{\hat{e}}_{V}{\hat{i}}_{V}}{4}{\cos \left( {{2\omega_{1}t} + \phi} \right)}}}} & (9) \end{matrix}$

where î_(diff0) is a DC current circulating through the two series-connected arms and the DC supply

$\begin{matrix} {{\hat{i}}_{{diff}\; 0} = \frac{{\hat{e}}_{V}{\hat{i}}_{V}\cos \; \phi}{u_{D} + \sqrt{u_{D}^{2} - {4R{\hat{\; e}}_{V}{\hat{i}}_{V}\cos \; \phi}}}} & (10) \end{matrix}$

When there is only a DC circulating current î_(diff0) then also the control voltage u_(diff)(t) becomes a DC voltage with the value u_(diff)(t)=Rî_(diff0) so that (3) becomes

$\begin{matrix} {{u_{CU}(t)} = {{\frac{u_{D}}{2} - {e_{V}(t)} - {R\; {\hat{i}}_{{{diff}\; 0}\mspace{14mu}}{u_{CL}(t)}}} = {\frac{u_{D}}{2} + {e_{V}(t)} - {R\; {\hat{i}}_{{diff}\; 0}}}}} & (11) \end{matrix}$

Moreover, equations (8) and (9) can be integrated, each with a freely selected integration constant, so that

$\begin{matrix} {{W_{CU}^{\Sigma}(t)} = {W_{{CU}\; 0}^{\Sigma} - {\frac{{\hat{e}}_{v}{\hat{i}}_{{diff}\; 0}}{\omega_{1}}\sin \; \omega_{1}t} + {\left( {\frac{u_{D}}{2} - {R\; {\hat{i}}_{{diff}\; 0}}} \right)\frac{{\hat{i}}_{V}}{2\omega_{1}}{\sin \left( {{\omega_{1}t} + \phi} \right)}} - {\frac{{\hat{e}}_{V}{\hat{i}}_{V}}{8\omega_{1}}{\sin \left( {{2\omega_{1}t} + \phi} \right)}}}} & (12) \\ {{W_{CL}^{\Sigma}(t)} = {W_{{CL}\; 0}^{\Sigma} + {\frac{{\hat{e}}_{v}{\hat{i}}_{{diff}\; 0}}{\omega_{1}}\sin \; \omega_{1}t} - {\left( {\frac{u_{D}}{2} - {R\; {\hat{i}}_{{diff}\; 0}}} \right)\frac{{\hat{i}}_{V}}{2\omega_{1}}{\sin \left( {{\omega_{1}t} + \phi} \right)}} - {\frac{{\hat{e}}_{V}{\hat{i}}_{V}}{4\omega_{1}}{\sin \left( {{2\omega_{1}t} + \phi} \right)}}}} & (13) \end{matrix}$

Thus the instantaneous energies in each arm can be calculated in real-time knowing only the references for the inner converter voltage and the actual AC current. The integration constants are the references for the desired average energy in each arm in the phase leg.

But if the energies are known then also the total capacitor voltage in the arms are know due to the connection equations

$\begin{matrix} {{u_{CU}^{\Sigma}(t)} = {{\sqrt{\frac{2{W_{CU}^{\Sigma}(t)}}{C_{arm}}}\mspace{14mu} {u_{CL}^{\Sigma}(t)}} = \sqrt{\frac{2{W_{CL}^{\Sigma}(t)}}{C_{arm}}}}} & (14) \end{matrix}$

Now the insertion indices valid for the desired steady-state operation can be calculated using equation (4). Given these insertion indices the energies in the upper and lower arms converges to the reference values given as free integration constants in (12) and (13). Normally these values are selected equal for both arms so that balanced operation is obtained. The value of the energy reference is selected to give the desired total capacitor voltage in each converter arm.

The second approach to stabilisation of the converter according to the procedure presented in this section is described in more detail in Appendix 2.

Remark 1: If a third harmonic voltage component (to increase the available voltage in a 3-phase converter) is added the formulas for the energies in the upper and lower arm will change somewhat. However the principle described in this paper still can be applied.

Remark 2: If even order harmonics are intentionally added to the circulating current the formulas for the energies in the upper and lower arms will change somewhat. The principle described above however still applies.

-   -   to derive a control strategy that provides main circuit         stability     -   to produce as high AC output voltage as possible with very low         harmonic distortion     -   to control the DC voltages of the capacitors in the modules

Continuous Model

It is of course possible to investigate the M2C converter by simulation. This approach however seems to be quite cumbersome in the sense that it involves detailed models of the arms (with tens of semiconductor devices in each). A lot of data will be generated making it more difficult to extract useful results.

Another approach, which will be followed here, makes use of a modulation principle that has been proposed by in DE10103031, in which a selection mechanism is used to determine which individual module that shall be inserted or bypassed when the number of devices in an arm shall be changed. The selection is made in dependence of the direction of the arm current (or phase current) and a comparison of the DC voltages in the congregation of modules in each arm, from which the modules having the highest and lowest voltages are identified.

Simulation has shown that this mechanism successfully keeps the DC voltages of the module capacitors quite close to each other, even for low number of modules in each arm (say e.g. five per arm). This functionality seems to remain even if the total switching frequency is low (a few hundred switchings per second for each semiconductor device).

Now it is assumed that this mechanism is in use and that accordingly there is no need to look at the DC voltages in the individual modules any more. The modulation process then can be described in terms of the total collective energy in each arm. As the total switching frequency (for all modules in each arm) becomes quite high continuous modelling can be used. The continuous model is a lot simpler to grasp than the detailed model and it is an ideal base for understanding the principles for the function of the M2C converter and to formulate control laws for different control aspects.

Due to the assumptions static relations exist between the total capacitor energy in the upper and lower arm, W_(CU) ^(Σ) and W_(CL) ^(Σ), and the corresponding total voltage of all capacitor modules in the arm, u_(CU) ^(Σ) and u_(CL) ^(Σ). Namely, if it is assumed that the energy is evenly shared between the modules, this relation becomes

$\begin{matrix} {{W_{CU}^{\Sigma} = {{N\left( {\frac{C}{2}\left( \frac{u_{CU}^{\Sigma}}{N} \right)^{2}} \right)} = {{\frac{C}{2N}u_{CU}^{\Sigma^{2}}\mspace{25mu} u_{CU}^{\Sigma}} = \sqrt{\frac{2N}{C}W_{CU}^{\Sigma}}}}}{W_{CL}^{\Sigma} = {{N\left( {\frac{C}{2}\left( \frac{u_{CL}^{\Sigma}}{N} \right)^{2}} \right)} = {{\frac{C}{2N}u_{CL}^{\Sigma^{2}}\mspace{25mu} u_{CL}^{\Sigma}} = \sqrt{\frac{2N}{C}W_{CL}^{\Sigma}}}}}} & \left( {A\; 1} \right) \end{matrix}$

where N is the number of modules per arm and C is the capacitance per module. In the following we will use the quantity ‘arm capacitance’ C_(arm) defined as follows

$\begin{matrix} {C_{arm} = \frac{C}{N}} & \left( {A\; 2} \right) \end{matrix}$

Then

$\begin{matrix} {{W_{CU}^{\Sigma} = {{\frac{C_{arm}}{2}\left( u_{CU}^{\Sigma} \right)^{2}\mspace{14mu} u_{CU}^{\Sigma}} = \sqrt{\frac{2W_{CU}^{\Sigma}}{C_{arm}}}}}W_{CL}^{\Sigma} = {\left. {\frac{C_{arm}}{2}\left( u_{CL}^{\Sigma} \right)^{2}}\mspace{11mu}\Leftrightarrow\; u_{CL}^{\Sigma} \right. = \sqrt{\frac{2W_{CL}^{\Sigma}}{C_{arm}}}}} & ({A3}) \end{matrix}$

Derivation of the Continuous Model

The electrical circuit representing the phase leg of the M2C converter is depicted in FIG. 2. The inserted capacitor voltages, U_(CU) and U_(CL), are created from the total capacitor voltages, u_(CU) ^(Σ) and u_(CL) ^(Σ), respectively, by applying the insertion indices, n_(U) and n_(L), which are controlled by the control system.

$\begin{matrix} {{n_{U} = {{\frac{u_{CU}}{u_{CU}^{\Sigma}}\mspace{14mu} 0} \leq n_{U} \leq 1}}{n_{L} = {{\frac{u_{CL}}{u_{CL}^{\Sigma}}\mspace{14mu} 0} \leq n_{L} \leq 1}}} & \left( {A\; 4} \right) \end{matrix}$

In the following, however, the main circuit model will be formulated using the real voltages as variables. If the total capacitor voltages, u_(CU) ^(Σ) and u_(CL) ^(Σ), are measured, the corresponding insertion indices can always be obtained from (A4).

The capacitor modules serve as controlled electromotive forces in the circuit. Let the total energy in the capacitors be W_(CU) ^(Σ) and W_(CL) ^(Σ) in the upper and lower arms respectively. Inspection of the circuit model in FIG. 2 immediately yields

$\begin{matrix} {{\frac{W_{CU}^{\Sigma}}{t} = {i_{U}u_{CU}}}{\frac{W_{CL}^{\Sigma}}{t} = {{- i_{L}}u_{CL}}}} & \left( {A\; 5} \right) \end{matrix}$

In order to gain some more insight it is helpful separate the arms currents in two parts. One part emerges from the AC current, which naturally separates into two halves, one passing though the upper and one passing through the lower arm. The deviation from this “ideal” condition is described by a difference current i_(diff) which passes through the series-connected arms and the DC source.

Define

$\begin{matrix} {{i_{U} = {\frac{i_{V}}{2} + i_{diff}}}{i_{L} = {\left. {\frac{i_{V}}{2} - i_{diff}}\Leftrightarrow i_{diff} \right. = \frac{i_{U} - i_{L}}{2}}}} & \left( {A\; 6} \right) \end{matrix}$

The circuit in FIG. 2 now gives the equations

$\begin{matrix} {{\frac{u_{D}}{2} - {R\left( {\frac{i_{V}}{2} + i_{diff}} \right)} - {L\left( {{\frac{1}{2}\frac{i_{V}}{t}} + \frac{i_{diff}}{t}} \right)} - u_{CU} - u_{V}} = {{0 - \frac{u_{D}}{2} - {R\left( {\frac{i_{V}}{2} - i_{diff}} \right)} - {L\left( {{\frac{1}{2}\frac{i_{V}}{t}} - \frac{i_{diff}}{t}} \right)} + u_{CL} - u_{V}} = 0}} & \left( {A\; 7} \right) \end{matrix}$

Adding and subtracting the equations give the results

$\begin{matrix} {{u_{V} = {\frac{u_{CL} - u_{CU}}{2} - {\frac{R}{2}i_{V}} - {\frac{L}{2}\frac{i_{V}}{t}}}}{{{L\frac{i_{diff}}{t}} + {Ri}_{diff}} = {\frac{u_{D}}{2} - \frac{u_{CL} + u_{CU}}{2}}}} & \left( {A\; 8} \right) \end{matrix}$

These equations show that

-   -   the AC voltage only depends on the AC current i_(V) and the         difference between the arm voltages u_(CL) and u_(CU)     -   the arm voltage difference acts as an inner AC voltage in the         converter and the inductance L and resistance R form a fix,         passive inner impedance for the AC current     -   the difference current i_(diff) only depends on the DC link         voltage and the sum of the arm voltages     -   the difference current i_(diff) can be controlled independently         of the AC side quantities by subtracting the same voltage         contributions to both arms

Define

$\begin{matrix} {{e_{V} = {{\frac{u_{CL} - u_{CU}}{2}\mspace{31mu} u_{CU}} = {\frac{u_{D}}{2} - e_{V} - u_{diff}}}}{u_{diff} = {\left. \frac{u_{D} - u_{CL} - u_{CU}}{2}\Leftrightarrow u_{CL} \right. = {\frac{u_{D}}{2} + e_{V} - u_{diff}}}}} & \left( {A\; 9} \right) \end{matrix}$

where e_(V) is the desired inner voltage in the AC voltage source and u_(diff) is a voltage that controls the difference current i_(diff).

Then (8) becomes

$\begin{matrix} {{u_{V} = {e_{V} - {\frac{R}{2}i_{V}} - {\frac{L}{2}\frac{i_{V}}{t}}}}{{{L\frac{i_{diff}}{t}} + {Ri}_{diff}} = u_{diff}}} & \left( {A\; 10} \right) \end{matrix}$

Inserting equations (A6) and (A9) in (A5) yields

$\begin{matrix} {{\frac{W_{CU}^{\Sigma}}{t} = {\left( {\frac{i_{V}}{2} + i_{diff}} \right)\left( {\frac{u_{D}}{2} - e_{V} - u_{diff}} \right)}}{\frac{W_{CL}^{\Sigma}}{t} = {\left( {{- \frac{i_{V}}{2}} + i_{diff}} \right)\left( {\frac{u_{D}}{2} + e_{V} - u_{diff}} \right)}}} & \left( {A\; 11} \right) \end{matrix}$

It makes sense to investigate the total energy stored in all capacitor banks in the whole leg and to examine the balance between the energy in the upper and the lower arm.

Define

$\begin{matrix} {{W_{C}^{\Sigma} = {{W_{CU}^{\Sigma} + {W_{CL}^{\Sigma}\mspace{31mu} W_{CU}^{\Sigma}}} = \frac{W_{C}^{\Sigma} + W_{C}^{\Delta}}{2}}}{W_{C}^{\Delta} = {\left. {W_{CU}^{\Sigma} - W_{CL}^{\Sigma}}\Leftrightarrow W_{CL}^{\Sigma} \right. = \frac{W_{C}^{\Sigma} - W_{C}^{\Delta}}{2}}}} & \left( {A\; 12} \right) \end{matrix}$

The result is

$\begin{matrix} {{\frac{W_{C}^{\Sigma}}{t} = {{\left( {u_{D} - {2\; u_{diff}}} \right)i_{diff}} - {e_{V}i_{V}}}}{\frac{W_{C}^{\Delta}}{t} = {{{- 2}\; e_{V}i_{diff}} + {\left( {\frac{u_{D}}{2} - u_{diff}} \right)i_{V}}}}} & \left( {A\; 13} \right) \end{matrix}$

Equation (A13) indicates that the total energy in both arms as well as the energy balance between the upper and lower arms can be controlled primarily by i_(diff), which is in its turn controlled by u_(diff) through (A10).

The term in the upper equation in (A13) is recognized as the instantaneous power delivered to the AC side.

p_(V)=e_(V)i_(V)  (A14)

Steady State Solution

We shall look at the special case where the AC emf and current is given. Thus let

$\begin{matrix} {{e_{V} = {{\hat{e}}_{V}\cos \mspace{11mu} \omega \; t}}{i_{V} = {{\hat{i}}_{V}\mspace{11mu} {\cos \left( {{\omega \; t} + \phi} \right)}}}} & ({A15}) \end{matrix}$

Assume that there is a solution where the difference current i_(diff) is a pure dc component. Thus

i _(diff)(t)=î _(diff)  (A16)

Then, according to (A10)

u _(diff)(t)=Rî _(diff)  (A17)

The derivative of the total and difference energies the according to (A13) become

$\begin{matrix} {{\frac{W_{C}^{\Sigma}}{t} = {{\left( {u_{D} - {2\; R{\hat{i}}_{diff}}} \right){\hat{i}}_{diff}} - {\frac{{\hat{e}}_{V}{\hat{i}}_{V}}{2}\left\lbrack {{\cos \; \phi} + {\cos \left( {{2\omega \; t} + \phi} \right)}} \right\rbrack}}}{\frac{W_{C}^{\Delta}}{t} = {{{- 2}\; {\hat{e}}_{V}{\hat{i}}_{diff}\cos \; \omega \; t} + {\left( {\frac{u_{D}}{2} - {R{\hat{i}}_{diff}}} \right){\hat{i}}_{V}{\cos \left( {{\omega \; t} + \phi} \right)}}}}} & \left( {A\; 18} \right) \end{matrix}$

From (A18) some observations can immediately be made

-   -   the derivative of the total energy contains only a constant and         a component having double network frequency     -   the derivative of the difference energy only contains components         having network frequency

Steady-state condition requires that the constant component of the total energy derivative disappears so that

$\begin{matrix} {{{{\left( {u_{D} - {2\; R{\hat{i}}_{diff}}} \right){\hat{i}}_{diff}} - {\frac{{\hat{e}}_{V}{\hat{i}}_{V}}{2}\cos \; \phi}} = 0}{{\hat{i}}_{diff} = \frac{{\hat{e}}_{V}{\hat{i}}_{V}\cos \; \phi}{u_{D} + \sqrt{u_{D}^{2} - {4\; R{\hat{e}}_{V}{\hat{i}}_{V}\cos \; \phi}}}}} & \left( {A\; 19} \right) \end{matrix}$

With this difference current the remaining term becomes

$\begin{matrix} {\frac{W_{C}^{\Sigma}}{t} = {{- \frac{{\hat{e}}_{V}{\hat{i}}_{V}}{2}}{\cos \left( {{2\omega \; t} + \phi} \right)}}} & \left( {A\; 20} \right) \end{matrix}$

The steady-state average energy can be freely selected so that the total energy in steady-state becomes

$\begin{matrix} {{W_{C}^{\Sigma}(t)} = {W_{C\; 0}^{\Sigma} - {\frac{{\hat{e}}_{V}{\hat{i}}_{V}}{4\omega}{\sin \left( {{2\omega \; t} + \phi} \right)}}}} & \left( {A\; 21} \right) \end{matrix}$

The expression for the difference energy In (A18) can be directly integrated, also with a freely selectable integration constant (which normally shall be zero)

$\begin{matrix} {{W_{C}^{\Delta}(t)} = {W_{C\; 0}^{\Delta} - {\frac{2{\hat{e}}_{V}{\hat{i}}_{diff}}{\omega}\sin \; \omega \; t} + {\frac{\left( {\frac{u_{D}}{2} - {R{\hat{i}}_{diff}}} \right){\hat{i}}_{V}}{\omega}{\sin \left( {{\omega \; t} + \phi} \right)}}}} & \left( {A\; 22} \right) \end{matrix}$

The investigation shows that

-   -   solutions of the desired type, i.e. with a difference component         having only a DC component, exist with freely selectable energy         levels in each arm     -   the steady state solutions for the energy time functions contain         only a double frequency component in the total energy and a         fundamental frequency component in the difference energy

Linearised Model for Control Studies

Let us go back and linearise the equations (A13) around a steady state point as described in the preceding section. Assume that the DC link voltage is constant. The

$\begin{matrix} {\mspace{79mu} {{\frac{{\Delta}\; W_{C}^{\Sigma}}{t} = {{\left( {u_{D} - {2\; R{\hat{i}}_{diff}}} \right)\Delta \; i_{diff}} - {2\; {\hat{i}}_{diff}\Delta \; u_{diff}} - {\Delta \; p_{V}}}}{\frac{{\Delta}\; W_{C}^{\Delta}}{t} = {{{- 2}e_{V}\Delta \; i_{diff}} - {i_{V}\Delta \; u_{diff}} - {2\; {\hat{i}}_{diff}\Delta \; e_{V}} + {\left( {\frac{u_{D}}{2} - {R{\hat{i}}_{diff}}} \right)\Delta \; i_{V}}}}}} & \left( {A\; 23} \right) \end{matrix}$

Further the differential equation that governs i_(diff) as function of u_(diff) applies so that

$\begin{matrix} {{{L\frac{{\Delta}\; i_{diff}}{t}} + {R\; \Delta \; i_{diff}}} = {\Delta \; u_{diff}}} & \left( {A\; 24} \right) \end{matrix}$

Stability Requirements

When the AC side current is stiff (A23) reduces to

$\begin{matrix} {{\frac{{\Delta}\; W_{C}^{\Sigma}}{t} = {{\left( {u_{D} - {2\; R{\hat{i}}_{diff}}} \right)\Delta \; i_{diff}} - {2\; {\hat{i}}_{diff}\Delta \; u_{diff}}}}{\frac{{\Delta}\; W_{C}^{\Delta}}{t} = {{{- 2}e_{V}\Delta \; i_{diff}} - {i_{V}\Delta \; u_{diff}}}}} & \left( {A\; 25} \right) \end{matrix}$

The linearized equations (A25) show that any control system, which makes the sum of the inserted voltages, u_(CL) and u_(CU), perfectly match the voltage u_(D) on the DC side, i.e. makes u_(diff)≡0, also makes the difference current become zero causing the derivatives of the energies in the arms to vanish. The main circuit in the converter then is marginally stable. Thus is not sufficient to select the inserted voltages in (A9) according to the desired e_(V), but an u_(diff) that creates stability must also be provided.

Control Law for the Total Capacitor Energy

The equation for the total energy equation can be formulated in the Laplace domain

$\begin{matrix} \begin{matrix} {{\Delta \; {W_{C}^{\Sigma}(s)}} = {{{\frac{1}{s}\left\{ {\frac{\left( {u_{D} - {2\; R{\hat{i}}_{diff}}} \right)}{R + {sL}} - {2\; {\hat{i}}_{diff}}} \right\} \Delta \; {u_{diff}(s)}} - \frac{\Delta \; {p_{V}(s)}}{s}} =}} \\ {= {{\frac{u_{D} - {4\; R{\hat{i}}_{diff}} - {2\; L{\hat{i}}_{diff}s}}{\left( {R + {sL}} \right)s}\Delta \; {u_{diff}(s)}} - \frac{\Delta \; {p_{V}(s)}}{s}}} \end{matrix} & \left( {A\; 26} \right) \end{matrix}$

Applying a Proportional Gain in an Energy Controller

Δu _(diff)(s)=K _(P) {ΔW _(C) ^(Σref)(s)−ΔW _(C) ^(Σ)(s)}  (A27)

yields

$\begin{matrix} {{\Delta \; {W_{C}^{\Sigma}(s)}} = {{\frac{\left\lbrack {u_{D} - {4\; R{\hat{i}}_{diff}} - {2\; L{\hat{i}}_{diff}s}} \right\rbrack K_{p}}{{\left( {R + {sL}} \right)s} + {\left\lbrack {u_{D} - {4\; R{\hat{i}}_{{diff}\;}} - {2\; L{\hat{i}}_{diff}s}} \right\rbrack K_{P}}}\Delta \; {W_{C}^{\Sigma \; {ref}}(s)}} - {\quad{{- \frac{R + {sL}}{{\left( {R + {sL}} \right)s} + {\left\lbrack {u_{D} - {4\; R{\hat{i}}_{diff}} - {2\; L{\hat{i}}_{diff}s}} \right\rbrack K_{P}}}}\Delta \; {p_{V}(s)}}}}} & \left( {A\; 28} \right) \end{matrix}$

The poles in the above transfer functions are mainly determined by

$\begin{matrix} {{{s_{1,2}^{2}L} + {u_{D}K_{p}}} = {\left. 0\Leftrightarrow s_{1,2} \right. = {{\pm j}\sqrt{\frac{u_{D}K_{p}}{L}}}}} & \left( {A\; 29} \right) \end{matrix}$

The control system is investigated for a converter leg in an example converter with the main parameters given in Table 1.

TABLE 1 Example converter main data 3-ph rated power 30 MVA rated frequency 50 Hz line-line voltage 13.8 kV rms rated phase current 1255 A rms arm capacitance 500 μF/arm arm inductance 3 mH arm resistance 100 mΩ

FIG. 3 shows the Nichols plot for the open loop transfer function in (A26) with the proportional gain K_(P)=0.001 V/J. The curve is almost independent of the active load.

As expected is the phase margin at 90 rad/s quite small, which means that the response will be quite oscillatory. In FIG. 4 the Nichols plot is shown when a PID controller is used.

The selected transfer function is given by

$\begin{matrix} {{F^{\Sigma}(s)} = {K_{P}\left( {1 + \frac{K_{I}}{s} + \frac{{sT}_{D}}{1 + {sT}_{F}}} \right)}} & \left( {A\; 30} \right) \end{matrix}$

with K_(P)=0.002 V/J, K_(I)=20 s⁻¹, T_(D)=10 ms, T_(F)=2 ms

It has been shown that the total energy response signal contains a frequency component with twice the network frequency. This component can be removed from the controller response using a notch filter. Further it is advisable to assume that a delay occurs in the measured total energy (total capacitor voltage). FIG. 5 shows the corresponding Nicols's diagram where the notch filter and the time delay have been included.

The transfer function in FIG. 5 is

$\begin{matrix} {{F^{\Sigma}(s)} = {{K_{P}\left( {1 + \frac{K_{I}}{s} + \frac{{sT}_{D}}{1 + {sT}_{F}}} \right)}^{- {sT}_{del}}\frac{s^{2} + \left( {2\omega_{1}} \right)^{2}}{s^{2} + {2{\zeta \left( {2\omega_{1}} \right)}s} + \left( {2\omega_{1}} \right)^{2}}}} & \left( {A\; 31} \right) \end{matrix}$

with K_(P)=0.002 V/J, K_(I)=20 s⁻¹, T_(D)=10 ms, T_(F)=2 ms, T_(del)=1 ms, ζ=0.05

FIG. 6 shows the simulation result at a step in the reference for the total energy in the converter leg.

Equation (A28) shows that the energy control system having only a proportional feedback will have a static error

$\begin{matrix} {\frac{\Delta \; W_{C}^{\Sigma}}{\Delta \; p_{V}} = \frac{- R}{K_{P}\left( {u_{D} - {4\; R{\hat{i}}_{diff}}} \right)}} & \left( {A\; 32} \right) \end{matrix}$

For the values in Table 1 together with K_(P)=0.002 V/J this energy dependence becomes approximately 0.002 J/W. Each leg of the converter handles about 10 MW causing the energy drop to be about 20 kJ (out of about 312 kJ) per leg.

FIG. 7 shows the simulation result when the current changes from 0.1 pu to 1.0 pu in the converter leg.

Control Law for Balancing the Capacitor Energies in the Arms

The general differential equation governing the balance between the energies in the upper and the lower arm was derived in (A13)

$\begin{matrix} {\frac{W_{C}^{\Delta}}{t} = {{{- 2}e_{V}i_{diff}} + {\left( {\frac{u_{D}}{2} - u_{diff}} \right)i_{V}}}} & \left( {A\; 33} \right) \end{matrix}$

It was linearised in (A23)

$\begin{matrix} {\frac{{\Delta}\; W_{C}^{\Delta}}{t} = {{{- 2}e_{V}\Delta \; i_{diff}} - {i_{V}\Delta \; u_{diff}} - {2{\hat{i}}_{diff}\Delta \; e_{V}} + {\left( {\frac{u_{D}}{2} - {R{\hat{i}}_{diff}}} \right)\Delta \; i_{V}}}} & \left( {A\; 34} \right) \end{matrix}$

If we consider linearising around the steady state solution defined by (A15) the linearised equation becomes

$\begin{matrix} {\frac{{\Delta}\; W_{C}^{\Delta}}{t} = {{{- 2}{\hat{e}}_{V}\cos \; \omega \; t\; \Delta \; i_{diff}} - {{\hat{i}}_{V}{\cos \left( {{\omega \; t} + \phi} \right)}\Delta \; u_{diff}} - {2{\hat{i}}_{diff}\Delta \; e_{V}} + {\left( {\frac{u_{D}}{2} - {R{\hat{i}}_{diff}}} \right)\Delta \; i_{V}}}} & \left( {A\; 35} \right) \end{matrix}$

Assume first that the AC side quantities are constant. Then

$\begin{matrix} {\frac{{\Delta}\; W_{C}^{\Delta}}{t} = {{{- 2}{\hat{e}}_{V}\cos \; \omega \; t\; \Delta \; i_{diff}} - {{\hat{i}}_{V}{\cos \left( {{\omega \; t} + \phi} \right)}\Delta \; u_{diff}}}} & \left( {A\; 36} \right) \end{matrix}$

Further assume that the controller produces a fundamental frequency sinusoidal signal with phase ζ relative the inner emf in the converter leg

u _(diff)(t)=û _(diff)(t)cos(ωt+ξ)  (A37)

Using the quasi-stationary solution to (A10) yields

$\begin{matrix} {{{{\hat{i}}_{diff}(t)} = {\frac{{\hat{u}}_{diff}}{\sqrt{R^{2} + \left( {\omega \; L} \right)^{2}}}{\cos \left( {{\omega \; t} + \xi - \eta} \right)}}}{\eta = {\arg \left( {R + {{j\omega}\; L}} \right)}}} & \left( {A\; 38} \right) \end{matrix}$

Inserting in (A36) we get

$\begin{matrix} {\frac{{\Delta}\; W_{C}^{\Delta}}{t} = {\left\{ {{{- \frac{2{\hat{e}}_{V}}{\sqrt{R^{2} + \left( {\omega \; L} \right)^{2}}}}{\cos \left( {\omega \; t} \right)}{\cos \left( {{\omega \; t} + \xi - \eta} \right)}} - {{\hat{i}}_{V}{\cos \left( {{\omega \; t} + \phi} \right)}{\cos \left( {{\omega \; t} + \xi} \right)}}} \right\} \Delta {\hat{u}}_{diff}}} & ({A39}) \end{matrix}$

The products of the cosine functions in (A39) are DC quantities and terms with the double network frequency. These components are

$\begin{matrix} {\mspace{79mu} {{a^{({d\; c})} = {{\frac{- {\hat{e}}_{V}}{\sqrt{R^{2} + \left( {\omega \; L} \right)^{2}}}{\cos \left( {\xi - \eta} \right)}} - {\frac{{\hat{i}}_{V}}{2}{\cos \left( {\xi - \phi} \right)}}}}{a^{({2\omega})} = {{\frac{- {\hat{e}}_{V}}{\sqrt{R^{2} + \left( {\omega \; L} \right)^{2}}}{\cos \left( {{2\omega \; t} + \xi - \eta} \right)}} - {\frac{{\hat{i}}_{V}}{2}{\cos \left( {{2\; \omega \; t} + \xi + \phi} \right)}}}}}} & ({A40}) \end{matrix}$

The relation between the two terms at various frequencies has been found to show that the first term dominates completely even for operating frequencies down to 5 Hz. Therefore it is sufficient to consider the first term. The maximum DC component then is obtained when

ξ=η=arg(R+jωL)  (A41)

With this selection of the argument for the inserted difference voltage we get the simplified formula

$\begin{matrix} {\frac{{\Delta}\; W_{C}^{\Delta}}{t} = {{- \frac{{\hat{e}}_{V}}{\sqrt{R^{2} + \left( {\omega \; L} \right)^{2}}}}\Delta \; {\hat{u}}_{diff}}} & ({A42}) \end{matrix}$

A proportional controller is sufficient to control the balance between the energies as the function is indeed just an integrator. However, the measured difference energy has a strong fundamental frequency component, which should be eliminated in the response to the regulator. Thus the transfer function in the balancing controller is given by

$\begin{matrix} {{F^{\Delta}(s)} = {K_{P\; \Delta}\frac{s^{2} + \omega_{1}^{2}}{s^{2} + {2{\zeta\omega}_{1}} + \omega_{1}^{2}}}} & ({A43}) \end{matrix}$

FIG. 8 shows the Nichols plot for the balance controller with parameters according to Table 1 and with control parameters K_(PΔ)=−0.005 V/J, t_(Del)=1 MS, ζ=0.1.

FIG. 5 shows that the closed loop for the energy controller has unity gain up to about 300 rad/s and that it amplifies frequencies in the range 100-200 rad/s with more than 3 dB. Therefore the gain in the balance controller has been kept low for these frequencies in order to avoid interaction between the two controllers.

Appendix 2: Description of Open Loop Control System

The aim of the investigation is to describe an M2C system where the modulation operates in open-loop mode. The meaning of the name “open-loop” in this context is that the modulation system does not measure the total voltage of the capacitors in the phase leg arms. Rather these total voltages are estimated in run-time using the desired AC emf and the measured AC current. The reference for the inserted arm voltages are obtained assuming that the instantaneous AC emf and AC current are steady state values. Further it is assumed that a voltage sharing system is provided to distribute the total arm voltage in each arm evenly between all modules that constitute the arm.

Steady State Analysis

The starting point is that the converter produces a sinusoidal emf

e_(V)=ê_(V) cos ω₁t  (B1)

and is loaded with a sinusoidal phase current

i _(V) =î _(V) cos(ω₁ t+φ)  (B2)

Under ideal conditions the arm currents only contains a DC component î_(diff0) so that the arm currents become

$\begin{matrix} {{i_{U} = {{\frac{{\hat{i}}_{V}}{2}{\cos \left( {{\omega_{1}t} + \phi} \right)}} + {\hat{i}}_{{diff}\; 0}}}{i_{L} = {{\frac{{\hat{i}}_{V}}{2}{\cos \left( {{\omega_{1}t} + \phi} \right)}} - {\hat{i}}_{{diff}\; 0}}}} & ({B3}) \end{matrix}$

When the difference current is î_(diff0) the difference voltage becomes u_(diff)=Rî_(diff0) so that the arm voltages become

$\begin{matrix} {{u_{CU} = {\frac{u_{D}}{2} - {{\hat{e}}_{V}\cos \; \omega_{1}t} - {R{\hat{i}}_{{diff}\; 0}}}}{u_{CL} = {\frac{u_{D}}{2} + {{\hat{e}}_{V}\cos \; \omega_{1}t} - {R{\hat{i}}_{{diff}\; 0}}}}} & ({B4}) \end{matrix}$

The derivatives of the arm energies are

$\begin{matrix} {{\frac{W_{CU}^{\Sigma}}{t} = {u_{CU}i_{U}}}{\frac{W_{CL}^{\Sigma}}{t} = {{- u_{CL}}i_{L}}}} & ({B5}) \end{matrix}$

Inserting the expressions in (B3) and (B4) yields

$\begin{matrix} {\frac{W_{CU}^{\Sigma}}{t} = {{{\left\{ {{\left( {\frac{u_{D}}{2} - {R{\hat{i}}_{{diff}\; 0}}} \right){\hat{i}}_{{diff}\; 0}} - \frac{{\hat{e}}_{V}{\hat{i}}_{V}\cos \; \phi}{4}} \right\}--}{\hat{e}}_{V}{\hat{i}}_{{diff}\; 0}\cos \; \omega_{1}t} + {\left( {\frac{u_{D}}{2} - {R{\hat{i}}_{{diff}\; 0}}} \right)\frac{{\hat{i}}_{V}}{2}{\cos \left( {{\omega_{1}t} + \phi} \right)}} - {\frac{{\hat{e}}_{V}{\hat{i}}_{V}}{4}{\cos \left( {{2\omega_{1}t} + \phi} \right)}}}} & ({B6}) \\ {\frac{W_{CL}^{\Sigma}}{t} = {\left\{ {{\left( {\frac{u_{D}}{2} - {R{\hat{i}}_{{diff}\; 0}}} \right){\hat{i}}_{{diff}\; 0}} - \frac{{\hat{e}}_{V}{\hat{i}}_{V}\cos \; \phi}{4}} \right\} - {{+ {\hat{e}}_{V}}{\hat{i}}_{{diff}\; 0}\cos \; \omega_{1}t} - {\left( {\frac{u_{D}}{2} - {R{\hat{i}}_{{diff}\; 0}}} \right)\frac{{\hat{i}}_{V}}{2}{\cos \left( {{\omega_{1}t} + \phi} \right)}} - {\frac{{\hat{e}}_{V}{\hat{i}}_{V}}{4}{\cos \left( {{2\omega_{1}t} + \phi} \right)}}}} & ({B7}) \end{matrix}$

In steady state the DC term must be zero. This condition allows us to determine the DC component to

$\begin{matrix} {{\hat{i}}_{{diff}\; 0} = \frac{{\hat{e}}_{V}{\hat{i}}_{V}\cos \; \phi}{u_{D} + \sqrt{u_{D}^{2} - {4R{\hat{e}}_{V}{\hat{i}}_{V}\cos \; \phi}}}} & ({B8}) \end{matrix}$

Thus, in steady state, the energy variations are

$\begin{matrix} {\frac{W_{CU}^{\Sigma}}{t} = {{{- {\hat{e}}_{V}}{\hat{i}}_{{diff}\; 0}\cos \; \omega_{1}t} + {\left( {\frac{u_{D}}{2} - {R{\hat{i}}_{{diff}\; 0}}} \right)\frac{{\hat{i}}_{V}}{2}{\cos \left( {{\omega_{1}t} + \phi} \right)}} - {\frac{{\hat{e}}_{V}{\hat{i}}_{V}}{4}{\cos \left( {{2\omega_{1}t} + \phi} \right)}}}} & ({B9}) \\ {\frac{W_{CL}^{\Sigma}}{t} = {{{+ {\hat{e}}_{V}}{\hat{i}}_{{diff}\; 0}\cos \; \omega_{1}t} - {\left( {\frac{u_{D}}{2} - {R{\hat{i}}_{{diff}\; 0}}} \right)\frac{{\hat{i}}_{V}}{2}{\cos \left( {{\omega_{1}t} + \phi} \right)}} - {\frac{{\hat{e}}_{V}{\hat{i}}_{V}}{4}{\cos \left( {{2\omega_{1}t} + \phi} \right)}}}} & ({B10}) \end{matrix}$

These formulas can immediately be integrated to obtain the instantaneous energy variations. Note that a freely selectable integration constant appears in each expression. Thus

$\begin{matrix} {{W_{CU}^{\Sigma}(t)} = {W_{{CU}\; 0}^{\Sigma} - {\frac{{\hat{e}}_{V}{\hat{i}}_{{diff}\; 0}}{\omega_{1}}\sin \; \omega_{1}t} + {\left( {\frac{u_{D}}{2} - {R{\hat{i}}_{{diff}\; 0}}} \right)\frac{{\hat{i}}_{V}}{2\omega_{1}}{\sin \left( {{\omega_{1}t} + \phi} \right)}} - {\frac{{\hat{e}}_{V}{\hat{i}}_{V}}{8\omega_{1}}{\sin \left( {{2\; \omega_{1}t} + \phi} \right)}}}} & ({B11}) \\ {{W_{CL}^{\Sigma}(t)} = {W_{{CL}\; 0}^{\Sigma} + {\frac{{\hat{e}}_{V}{\hat{i}}_{{diff}\; 0}}{\omega_{1}}\sin \; \omega_{1}t} - {\left( {\frac{u_{D}}{2} - {R{\hat{i}}_{{diff}\; 0}}} \right)\frac{{\hat{i}}_{V}}{2\omega_{1}}{\sin \left( {{\omega_{1}t} + \phi} \right)}} - {\frac{{\hat{e}}_{V}{\hat{i}}_{V}}{4\omega_{1}}{\sin \left( {{2\; \omega_{1}t} + \phi} \right)}}}} & ({B12}) \end{matrix}$

The total capacitor voltages now are given by

$\begin{matrix} {{{u_{CU}^{\Sigma}(t)} = \sqrt{\frac{2{W_{CU}^{\Sigma}(t)}}{C_{arm}}}}{{u_{CL}^{\Sigma}(t)} = \sqrt{\frac{2{W_{CL}^{\Sigma}(t)}}{C_{arm}}}}} & ({B13}) \end{matrix}$

and they can be used to determine the insertion indices in run-time according to

$\begin{matrix} {{{n_{U}(t)} = \frac{u_{CU}(t)}{u_{CU}^{\Sigma}(t)}}{{n_{L}(t)} = \frac{u_{CL}(t)}{u_{CL}^{\Sigma}(t)}}} & ({B14}) \end{matrix}$

Open-Loop Control

The idea of the open-loop control mode is to

-   -   measure the AC terminal current     -   extract the amplitude and phase relative the created emf in the         converter as in (B2)     -   perform the calculation as described above     -   utilize the so obtained insertion indices according to (B14) in         the converter

The invention has mainly been described above with reference to a few embodiments. However, as is readily appreciated by a person skilled in the art, other embodiments than the ones disclosed above are equally possible within the scope of the invention, as defined by the appended patent claims. 

1. A method for calculating insertion indices for a phase leg of a DC to AC modular multilevel converter, the converter comprising one phase leg between upper and lower DC source common bars for each phase, each phase leg comprising two serially connected arms, wherein an AC output for each phase leg is connected between its two serially connected arms, wherein each arm comprises a number of submodules, wherein each submodule can be in a bypass state or a voltage insert mode, the insertion index comprising data representing the portion of available submodules that should be in the voltage insert mode for a particular arm, the method comprising the steps of: calculating a desired arm voltage for an upper arm connected to the upper DC source common bar and a lower arm connected to the lower DC source common bar, obtaining values representing actual total arm voltages in the upper arm and lower arm, respectively, and calculating insertion indices for the upper and lower arm, respectively, using the respective desired arm voltage and the respective value representing the total actual arm voltage; wherein the step of calculating desired arm voltages for a phase leg comprises calculating u _(CU)(t)=u _(D)/2−e _(V)(t)−u _(diff)(t) for the upper arm, and calculating u _(CL)(t)=u _(D)/2+e _(V)(t)−u _(diff)(t) for the lower arm, where u_(CU)(t) represents desired upper arm voltage where u_(CL)(t) represents desired lower arm voltage, u_(D) represents a voltage between the upper and lower DC source common bars, e_(V)(t) represents a reference inner AC output voltage and u_(diff)(t) represents a control voltage to control a current passing through the whole phase leg, and calculating u _(diff)(t)=u _(diff1)(t)+u _(diff2)(t) where u_(diff1)(t) represents a voltage obtained by summing energy in the arms of the leg and u_(diff2)(t) represents a voltage obtained by calculating a difference in energy between the arms of the leg.
 2. The method according to claim 1, wherein the step of obtaining a value representing actual arm voltage comprises calculating u _(diff2)(t)=û _(diff2) cos(ω₁ t+ψ) where û_(diff2) represents a difference between total upper arm energy and total lower arm energy, ω₁ represents the angular velocity of the network frequency and ψ represents the angle given by ψ=<(R+jω₁L) where R represents the resistance of the converter arm and L represents the inductance of the converter arm.
 3. A method for calculating insertion indices for a phase leg of a DC to AC modular multilevel converter, the converter comprising one phase leg between upper and lower DC source common bars for each phase, each phase leg comprising two serially connected arms, wherein an AC output for each phase leg is connected between its two serially connected arms, wherein each arm comprises a number of submodules, wherein each submodule can be in a bypass state or a voltage insert mode, the insertion index comprising data representing the portion of available submodules that should be in the voltage insert mode for a particular arm, the method comprising the steps of: calculating a desired arm voltage for an upper arm connected to the upper DC source common bar and a lower arm connected to the lower DC source common bar, obtaining values representing actual total arm voltages in the upper arm and lower arm, respectively, and calculating insertion indices for the upper and lower arm, respectively, using the respective desired arm voltage and the respective value representing the total actual arm voltage; wherein the step of calculating desired arm voltages for a phase leg comprises calculating u _(CU)(t)=u _(D)/2−e _(V)(t)−u _(diff)(t) for the upper arm, and calculating u _(CL)(t)=u _(D)/2+e _(V)(t)−u _(diff)(t) for the lower arm, where u_(CU)(t) represents desired upper arm voltage where u_(CL)(t) represents desired lower arm voltage, u_(D) represents a voltage between the upper and lower DC source common bars, e_(V)(t) represents a reference inner AC output voltage and u_(diff)(t) represents a control voltage to control a current passing through the whole phase leg, wherein the step of obtaining values representing actual arm voltages comprises: calculating u_(CU) ^(Σ)(t), actual total voltage for the upper arm, using C_(arm), capacitance for the arm, î_(diff0), DC current passing through the two serially connected arms of the phase leg, W_(CU) ^(Σ)(t), desired average energy in the upper arm, ê_(V), amplitude of reference for the inner AC output voltage, î_(V), amplitude of AC output current, φ, a phase difference between i_(V)(t) and e_(V)(t), DC current circulating through the two series-connected arms, and calculating u_(CL) ^(Σ)(t), actual total voltage for the lower arm, using C_(arm), capacitance for the arm, î_(diff0), DC current passing through the two serially connected arms of the phase leg, W_(CL) ^(Σ)(t), desired average energy in the lower arm, ê_(V), amplitude of reference for inner AC output voltage, î_(V), amplitude of AC output current φ, a phase difference between i_(V)(t) and e_(V)(t), î_(diff0), DC current circulating through the two series-connected arms.
 4. The method according to claim 3, wherein the step of obtaining a value representing actual arm voltage comprises calculating ${\hat{i}}_{{diff}\; 0} = \frac{{\hat{e}}_{v}{\hat{i}}_{v}\cos \; \phi}{u_{D} + \sqrt{u_{D}^{2} - {4\mspace{11mu} R{\hat{e}}_{v}{\hat{i}}_{v}\cos \; \phi}}}$ where φ represents a phase difference between i_(V)(t) and e_(V)(t), u_(D) represents a voltage between the upper and lower DC source common bars and R represents the resistance of the converter arm.
 5. The method according to claim 1, wherein the step of obtaining a value representing actual total arm voltage comprises measuring voltages of the submodules of the arm and summing these measured voltages.
 6. The method according to claim 1, wherein the insertion index comprises data representing a direction of the inserted voltage.
 7. An apparatus for calculating insertion indices for a phase leg of a DC to AC modular multilevel converter, the converter comprising one phase leg between upper and lower DC source common bars for each phase, each phase leg comprising two serially connected arms, wherein an AC output for each phase leg is connected between its two serially connected arms, wherein each arm comprises a number of submodules, wherein each submodule can be in a bypass state or a voltage insert mode, the insertion index comprising data representing the portion of available submodules that should be in the voltage insert mode for a particular arm, the apparatus comprises: a controller arranged to calculate a desired arm voltage for an upper arm connected to the upper DC source common bar and a lower arm connected to the lower DC source common bar, to obtain values representing actual total arm voltages in the upper arm and lower arm, respectively, and to calculate insertion indices for the upper and lower arm, respectively, using the respective desired arm voltage and the respective value representing the total actual arm voltage; wherein the calculating of desired arm voltages for a phase leg comprises calculating u _(CU)(t)=u _(D)/2−e _(V)(t)−u _(diff)(t) for the upper arm, and calculating u _(CL)(t)=u _(D)/2+e _(V)(t)−u _(diff)(t) for the lower arm, where u_(CU)(t) represents desired upper arm voltage, u_(CL)(t) represents desired lower arm voltage, u_(D) represents a voltage between the upper and lower DC source common bars, e_(V)(t) represents a reference inner AC output voltage and u_(diff)(t) represents a control voltage to control a current passing through the whole phase leg, and calculating u _(diff)(t)=u _(diff1)(t)+u _(diff2)(t) where u_(diff1)(t) represents a voltage obtained by summing energy in the arms of the leg and u_(diff2)(t) represents a voltage obtained by calculating a difference in energy between the arms of the leg.
 8. An apparatus for calculating insertion indices for a phase leg of a DC to AC modular multilevel converter, the converter comprising one phase leg between upper and lower DC source common bars for each phase, each phase leg comprising two serially connected arms, wherein an AC output for each phase leg is connected between its two serially connected arms, wherein each arm comprises a number of submodules, wherein each submodule can be in a bypass state or a voltage insert mode, the insertion index comprising data representing the portion of available submodules that should be in the voltage insert mode for a particular arm, the apparatus comprises: a controller arranged to calculate a desired arm voltage for an upper arm connected to the upper DC source common bar and a lower arm connected to the lower DC source common bar, to obtain values representing actual total arm voltages in the upper arm and lower arm, respectively, and to calculate insertion indices for the upper and lower arm, respectively, using the respective desired arm voltage and the respective value representing the total actual arm voltage; wherein the calculating of desired arm voltages for a phase leg comprises calculating u _(CU)(t)=u _(D)/2−e _(V)(t)−u _(diff)(t) for the upper arm, and calculating u _(CL)(t)=u _(D)/2+e _(V)(t)−u _(diff)(t) for the lower arm, where u_(CU)(t) represents desired upper arm voltage, u_(CL)(t) represents desired lower arm voltage, u_(D) represents a voltage between the upper and lower DC source common bars, e_(V)(t) represents a reference inner AC output voltage and u_(diff)(t) represents a control voltage to control a current passing through the whole phase leg, wherein the step of obtaining values representing actual arm voltages comprises: calculating u_(CU) ^(Σ)(t), actual total voltage for the upper arm, using C_(arm), capacitance for the arm, î_(diff0), DC current passing through the two serially connected arms of the phase leg, W_(CU) ^(Σ)(t), desired average energy in the upper arm, ê_(V), amplitude of reference for the inner AC output voltage, î_(V), amplitude of AC output current, φ, a phase difference between i_(V)(t) and e_(V)(t), DC current circulating through the two series-connected arms, and calculating u_(CL) ^(Σ)(t), actual total voltage for the lower arm, using C_(arm), capacitance for the arm, î_(diff0), DC current passing through the two serially connected arms of the phase leg, W_(CL) ^(Σ)(t), desired average energy in the lower arm, ê_(V), amplitude of reference for inner AC output voltage, î_(V), amplitude of AC output current φ, a phase difference between i_(V)(t) and e_(V)(t), î_(diff0), DC current circulating through the two series-connected arms. 